ENGENHARIA DE MATERIAIS
SPRAY FORMING OF WEAR AND CORROSION RESISTANT BIMETALLIC PIPES: FROM THE ALLOY DESIGN TO THE SEMI-INDUSTRIAL PROCESS
Guilherme Zepon
UNIVERSIDADE FEREDAL DE SÃO CARLOS CENTRO DE CIÊNCIAS EXATAS E DE TECNOLOGIA
PROGRAMA DE PÓS-GRADUAÇÃO EM CIÊNCIA E ENGENHARIA DE MATERIAIS
SPRAY FORMING OF WEAR AND CORROSION RESISTANT BIMETALLIC PIPES: FROM THE ALLOY DESIGN TO THE SEMI-INDUSTRIAL PROCESS
Guilherme Zepon
Tese apresentada ao Programa de Pós-Graduação em Ciência e Engenharia de Materiais como requisito parcial à obtenção do título de DOUTOR EM CIÊNCIA E ENGENHARIA DE MATERIAIS
Orientador: Dr. Claudemiro Bolfarini Agência Financiadora: FAPESP
Z57s Spray forming of wear and corrosion resistant bimetallic pipes : from the alloy design to the semi-industrial process / Guilherme Zepon. -- São Carlos : UFSCar, 2016.
169 p.
Tese (Doutorado) -- Universidade Federal de São Carlos, 2016.
DEDICATION
This doctoral thesis is dedicated to my family, Eleodora, Antonio Carlos, Muriel Tamires and Cibele.
VITAE DO CANDIDATO
AKNOWLEDGEMENT
I would like to thank FAPESP for its financial support. Doctorate scholarship (Grant number: 12/25352-4) and the scholarship abroad (Grant number: 14/07384-1).
I am very grateful to my supervisor Professor Claudemiro Bolfarini for his advice, support and encouragement.
I would like to thank Dr. Volker Uhlenwinkel and Prof. Ricardo Nogueira for welcoming me in their laboratories and for all attention they have given me during my period abroad.
I also would like to thank the laboratory technicians Beto and Lemão for all support in the practical work.
Special thanks to my friends of our research group Eric, Banzo, Brunão, Braulio, Otani, Pama, Ferrugem, Witor, Negão and the LCE staff.
I am also deeply grateful to my dear friends Mono, Bia, Zigoto, Fer, Jeremias, Márcio, Tuco and Jana for sharing with me special moments.
ABSTRACT
The oil exploitation and production at the pre-salt fields in a safety and efficient way depends on the development of materials that withstand the severe work conditions found in these fields. For instance, pipes, such as drilling risers and casings, are often subjected to severe wear and corrosion conditions. This thesis is dedicated to evaluate the technical feasibility to produce wear and corrosion resistant bimetallic pipes by spray forming. The processing-microstructure-properties relationship of the spray-formed boron-modified supermartensitic stainless steel (SMSS) grades was comprehensively studied. Deposits of SMSS with boron contents ranging from 0.3 %wt. to 1.0 %wt. were processed by spray forming. The spray-formed boron-modified SMSS deposits had the wear resistance evaluated through different wear tests and their corrosion resistances by means of electrochemical techniques. It was demonstrated that the wear resistance of the spray-formed boron-modified SMSS is determined by the presence of the eutectic network of M2B-type borides resulted from the spray
CONFORMAÇÃO POR SPRAY DE TUBOS BIMETÁLICOS RESISTENTES AO DESGASTE E À CORROSÃO: DA CONCEPÇÃO DAS LIGAS AO
PROCESSAMENTO SEMI-INDUSTRIAL RESUMO
A exploração e produção de petróleo nos poços do pré-sal de modo seguro e eficiente dependem do desenvolvimento de ligas que suportem severas condições de trabalho. Por exemplo, tubos como risers de perfuração e casings são frequentemente submetidos à severas condições de desgaste e corrosão. Esta tese se dedica a avaliar a viabilidade técnica de produzir tubos bimetálicos resistentes à corrosão e ao desgaste conformados por spray. A relação processamento-microestrutura-propriedade do aço inoxidável supermartensítico (AISM) modificado com boro e conformado por spray foi estudada de forma abrangente. Depósitos de AISM com teores de boro variando de 0,3%p. a 1,0%p. foram processados por conformação por spray. Os depósitos tiveram a resistência ao desgaste avaliada através de diferentes ensaios e a resistência à corrosão por meio de técnicas eletroquímicas. Demonstrou-se que a resistência ao desgaste das ligas modificadas com boro é determinada pela presença de boretos eutéticos, do tipo M2B, oriunda do processo de conformação por spray.
PUBLICATIONS
- ZEPON, G.; NASCIMENTO, A.R.C.; KASAMA, A.H.; NOGUEIRA, R.P.; KIMINAMI, C.S.; BOTTA, W.J.; BOLFARINI, C. Design of wear resistant boron-modified supermartensitic stainless steel by spray forming process. Materials & Design 83 (2015) 214-223.
- ZEPON, G.; ELLENDT, N.; UHLENWINKEL, V.; BOLFARINI, C. Solidification sequence of spray-formed steels. Metallurgical and Materials Transactions A 47 (2016) 842-851.
INDEX
FOLHA DE APROVAÇÃO... i
AKNOWLEDGEMENT ... iii
ABSTRACT ... v
RESUMO ... vii
PUBLICATIONS ... ix
TABLE INDEX ... xiii
FIGURE INDEX ... xv
1 INTRODUCTION ... 1
2 LITERATURE ... 5
2.1 Spray Forming and Co-Spray Forming Fundamentals ... 5
2.2 Supermartensitic Stainless Steel ... 17
2.3 Boron modified stainless steels ... 22
3 OBJECTIVES ... 27
4 MATERIALS AND METHODS ... 29
4.1 Design of Spray-formed boron-modified SMSS ... 29
4.1.1 Laboratory Scale Spray Forming ... 29
4.1.2 Microstructure Characterization ... 31
4.1.3 Equilibrium Solidification path determination ... 32
4.1.4 Wear tests ... 32
4.1.5 Electrochemical corrosion tests ... 35
4.2 Spray forming of semi-industrial scale bimetallic pipes ... 38
4.2.1 Spray Forming runs ... 38
4.2.2 Porosity and microstructure characterization ... 43
4.3.1 Hardness and Heat treatments ... 44
4.3.2 Sampling and mechanical tests ... 44
5 RESULTS AND DISCUSSION ... 47
5.1 Design of spray-formed boron modified SMSS ... 47
5.1.1 Deposits characterization ... 47
5.1.2 Microstructural evolution in spray forming ... 53
5.1.3 Effect of boron content on the wear resistance ... 64
5.1.3.1Dry sand against rubber wheel abrasive wear test ... 64
5.1.3.2 Plate-on-cylinder wear test ... 67
5.1.4 Effect of boron content on the corrosion Resistance ... 73
5.2 Spray Forming of semi-industrial bimetallic pipes ... 87
5.2.1 Effect of process parameters on the shape, porosity, and surface temperature ... 87
5.2.2 Effect of temperature profile on the pipe microstructure. ... 101
5.2.3 Effect of temperature profile on the interface’s porosity. ... 111
5.3 Mechanical Properties ... 119
5.3.1 Hardness and Heat treatments ... 119
5.3.2 Mechanical properties of the SMSS layer ... 125
5.3.3 Mechanical properties of the SM-14Cr-1B layer ... 133
6 GENERAL DISCUSSION ... 139
7 CONCLUSIONS ... 149
8 SUGGESTION FOR FUTURE WORKS ... 151
9 REFERENCES ... 153
TABLE INDEX
Table 2.1 - Materials properties and standard boundary conditions used for the numerical simulation of spray forming of CuSn6 and 100Cr6 steel billets [10,50]. ... 8 Table 2.2 - Typical chemical composition of SMSS grades [56]... 19 Table 2.3 - Alloy design of the SMSS grades to meet target corrosion resistance
[57]. ... 19 Table 2.4 - Mechanical properties of the three SMSS grades. Hot rolled plates
after optimized heat treatments with tempering temperatures close to AC1
[56]. ... 21
Table 4.1 - Aimed chemical composition of the four spray-formed boron-modified SMSS. ... 31 Table 4.2 - Composition of the drilling MUD used in the plate-on-cylinder wear
tests. ... 33 Table 4.3 - Chemical composition of the commercial supermartensitic stainless
steel (%wt.). ... 38 Table 4.4 - Chemical composition (%wt.) of the feedstock materials. ... 39 Table 4.5 - Amounts of feedstock material used in each spray-forming run. .... 40 Table 4.6 - Target chemical composition of the spray-formed alloys. ... 40 Table 4.7 - Process parameters of the semi-industrial scale spray-forming runs. ... 42
Table 5.1 - Chemical composition of the spray-formed boron-modified SMSSs. ... 48 Table 5.2 - EDS microanalyses of the martensitic matrixes and the M2B-type
borides of the spray-formed boron-modified SMSSs. ... 52 Table 5.3 - Fitted parameters from experimental POC wear test data. ... 68 Table 5.4 - Mean values of the corrosion properties obtained from the polarization
Table 5.5 - Tensile properties of the of the SMSS layer as-spray formed and tempered at 650 ºC for 2h. ... 127 Table 5.6 - Tensile properties of the as-spray formed and tempered at 650 ºC for
FIGURE INDEX
Figure 1.1 - (a) Bending of the drilling riser due to the movement of the drilling platform. (b) Wear of the drilling riser wall caused by the contact with the rotating tool joint (connectors of the drill pipe). ... 1 Figure 1.2 - (a) Wear grooves caused by the contact of rotating tool joints in the
inner wall of a casing. (b) Drilling riser failed by thickness loss caused by wear and corrosion. ... 2
Figure 2.1 - Schematic representation of spray forming process. ... 5 Figure 2.2 - Numerical simulation results showing the overall temperature at
different times of (a) CuSn6 and (b) 100Cr6 steel (AISI 52100) billets spray-formed using similar process conditions. Total time of the spraying process: 360 s [10,50]. ... 9 Figure 2.3 - Surface temperature distribution and temperature distribution through
the longitudinal section of tubular preform of 100Cr6 steel (AISI 52100) under standard spray condition at spraying time of: (a) and (d) 30 s; (b) and (e) 90 s; (c) and (f) 120 s [20]... 12 Figure 2.4 - Effect of the enthalpy input from the spray on the thermal profiles of
a 100Cr6 steel (AISI 52100) tubular preform at the deposition time of 120 s with average liquid fraction in the spray of (a) 0.3 and (b) 0.7 [20]. ... 12 Figure 2.5 - Porosity versus deposit surface temperature of spray-formed Ni
superalloy rings [45]. ... 14 Figure 2.6 - Effect of the dimensionless enthalpy of deposit surface on porosity
of Al-bronze, Sn-bronze and nitriding steel [16]. ... 14 Figure 2.7 - Schematic representation of co-spray forming process by-layer (a)
tubes and (b) plates. ... 15 Figure 2.8 - Cross sections of a Stellite21/AISI H13 steel co-spray formed tube
processed with different parameters and the results of their liquid penetrant inspection [53]. ... 16 Figure 2.9 - Schematic of co-spray forming of a gradient deposit from two different
Figure 2.10 - (a) Effect of molybdenum on CO2 corrosion resistance at 180 ºC for
two different nickel concentration. (b) Effect of molybdenum on the resistance to sulfide stress-corrosion cracking (applied stress: 100% yield strength). Base composition 0.025C-13Cr-0.45Mn (wt%.) [58]. ... 20 Figure 2.11 - Experimental diagram showing the boundaries of the austenite,
ferrite and martensite phases as function of Cr, Ni and Mo concentration for 0.01 wt.% C after austenitization at 1050 ºC and air cooling [55]. ... 20 Figure 2.12 - Impact resistance (Charpy-V) at sub-zero temperatures of the three
typical SMSS grades. Optimized microstructure free of δ-ferrite. The hot rolled plates were heated over 800 ºC and water quenched. The optimum balance between strength and toughness is achieved after tempering in a narrow range of temperatures close to AC1. Typical grain size is ASTM:
7-10. ... 22 Figure 2.13 - Microstructure of the spray-formed SDSS modified with 3.5%wt. of
boron [1]. ... 23 Figure 2.14 - Microstructure of the spray-formed AISI 430 modified with (a) 1.0
%wt., (b) 2.0 %wt. and (c) 4.0 %wt. of boron [4]. ... 24 Figure 2.15 - Microstructure of spray-formed SMSS modified with (a) 0.3 %wt.
and (b) 0.7%wt. of boron [7]. ... 25
Figure 4.1 - Close-coupled spray forming equipment of the DEMa-UFSCar. ... 30 Figure 4.2 - Schematic representation of the plate-on-cylinder wear test. ... 33 Figure 4.3 - Microstructure of the commercial SMSS bar used as reference alloy
for the wear tests. Etching: Vilela. ... 35 Figure 4.4 - (a) Schematic representation and (b) photography of the
electrochemical cell used for potentiodynamic and EIS analyses. ... 37 Figure 4.5 - Schematic representation of the co-spray forming process performed
at the SK1+ plant in the IWT showing the induction preheating system and optical pyrometer positioning. ... 38 Figure 4.6 - Schematic representation of the pyrometer measurement range
Figure 4.7 - Representation of sampling regions of the SF3 bimetallic pipe for mechanical tests. ... 45 Figure 4.8 - Dimensions of the sub-sized tension test specimen used in this work. ... 45
Figure 5.1 - Spray-formed boron-modified SMSS deposits produced at the close-coupled spray forming equipment at DEMa-UFScar. ... 47 Figure 5.2 - (a) XRD patterns of the spray-formed boron-modified SMSSs. (b)
Zoom of the SM-1B and SM-14Cr-1B XRD patterns. ... 49 Figure 5.3 - Microstructures and borides morphology of (a) (b) SM-0.3B, (c) (d)
SM-0.7B, (e) (f) SM-1B and (g) (h) SM-14Cr-1B observed by SEM images (secondary electrons)... 50 Figure 5.4 - Grain size and hardness of the spray-formed boron-modified SMSS. ... 52 Figure 5.5 - Pseudo-binary phase diagram of the boron-modified SMSS medium-alloyed grade (Fe-12Cr-5Ni-2Mo). ... 54 Figure 5.6 - Calculated amount of phases (in mole fraction) showing the
equilibrium solidification path of the (a) SM-0.3B, (b) SM-0.7B, (c) SM-1B, and (d) SM-14Cr-1B. ... 55 Figure 5.7 - Microstructure of the overspray powder of SM-1B with (a) 200 µm
(OM) and (b) 50 µm (SEM – secondary electrons image). ... 57 Figure 5.8 - Pseudo-binary phase diagram showing the solidification path of the
SM-1B. TL = temperature of the fully liquid droplets; TPL = temperature of
partially solidified droplets; TS = temperature of completely solidified droplets;
and Teq = equilibrium temperature of the deposition zone... 59
Figure 5.11 - SEM images (secondary electrons) of the worn surfaces of (a) commercial SM, (b) SM-0.3B, (c) SM-0.7B and (d) SM-1B after the dry sand against rubber wheel abrasive wear test. ... 65 Figure 5.12 - Schematic illustration of the three-body wear mechanism in the dry
sand against rubber wheel wear test of (a) SMSS, (b) 0.3B, and (c) SM-1B. ... 66 Figure 5.13 - Experimental data and fitted curves of accumulated worn volume
versus sliding distance in the POC wear test. ... 68 Figure 5.14 - Samples of the (a) commercial SMSS, (b) SM-0.3B, (c) SM-0.7B,
and (d) SM-1B after the POC wear test. ... 69 Figure 5.15 - (a) Wear rate and (b) contact pressure versus sliding distance in
the plate-on-cylinder wear test. ... 71 Figure 5.16 - SEM images (secondary electrons) of the worn surfaces and plate-like debris of (a) and (b) commercial SMSS, (c) and (d) SM-0.3B, (e) and (f) SM-0.7B and (g) and (h) SM-1B after the POC wear test. ... 72 Figure 5.17 - (a) EIS and (b) polarization results of the commercial SMSS.
Electrolyte: 35 g/L NaCl, pH = 4.0. ... 74 Figure 5.18 - (a) EIS and (b) polarization results of the spray-formed boron-modified SMSSs in protocol I. ... 77 Figure 5.19 - (a) EIS and (b) polarization results of the spray-formed boron-modified SMSSs in protocol II. ... 78 Figure 5.20 - (a) EIS and (b) polarization results of the spray-formed boron-modified SMSSs in protocol III. ... 80 Figure 5.21 - Corroded surfaces of (a) SM-0.3B, (b) SM-0.7B, (c) SM-1B and (d)
SM-14Cr-1B. ... 81 Figure 5.22 - Coupling currents measured by EN before and after abrasion of (a)
commercial SMSS, (b) SM-0.3B, and (b) SM-14Cr-1B. Arbitrary zero time origin. ... 83 Figure 5.23 - (a) I/Imax and (b) I/Icorr measured by EN after abrasion of the
Figure 5.24 - (a) Spray-formed SM-13Cr-1B/SMSS bimetallic pipe produced in the SF1 run. (b) Transversal section of the bimetallic section. ... 87 Figure 5.25 - Sequence showing the fracturing and expelling of the SM-13Cr-1B
layer during deposition in SF1. ... 88 Figure 5.26 - (a) Porosity profile of the bimetallic pipe spray formed in SF1 run.
(b) Zoom of the porosity profile of the SM-13Cr-1B layer... 89 Figure 5.27 - OM images showing the porosity profile of the SF1 spray-formed
pipe. (a) Middle of the SMSS layer, (b) interface region, (c) middle of the SM-13Cr-1B layer, (d) bottom of the SM-SM-13Cr-1B layer. (e) Zoom of the porosity in the SMSS layer. ... 90 Figure 5.28 - (a) Spray-formed SM-14Cr-1B pipe produced in the SF2 run. (b)
Transversal section at the middle of pipe (dashed lined in (a)). ... 92 Figure 5.29 - Porosity profile of the SM-14Cr-1B pipe spray-formed in the SF2
run. ... 92 Figure 5.30 - (a) Spray-formed SM-14Cr-1B/SMSS bimetallic pipe produced in
the SF3 run. (b) Transversal section at the middle of pipe (dashed lined in (a)). (c) Zoom detailing the porous zone at the bi-layer tube. ... 94 Figure 5.31 - Porosity profile of the bimetallic pipe spray-formed in the SF3 run. ... 95 Figure 5. 32 - OM images showing the porosity profile of the SF3 spray-formed
pipe. Region (a) near the top surface, (b) middle, and (c) near the interface of the SMSS layer. (d) Middle and (e) bottom region of the SM-14Cr-1B layer. ... 96 Figure 5.33 - Surface temperature charts from the optical pyrometer
measurements. (a) SF2 and (b) SF3 run. ... 99 Figure 5.34 - Surface temperature (a) at the center of the first spray cone -
Figure 5.36 - Microstructure and borides morphology (deep etching) of the SM-13Cr-1B layer from the SF1 run at (a) and (b) bottom part (1 mm from substrate); and (c) and (d) middle part (6 mm from substrate). ... 103 Figure 5.37 - OM images showing the microstructure of the SM-14Cr-1B SF2 pipe
at (a) 3 mm, (b) 6 mm, and (c) 10 mm from the substrate; and the SM-14Cr-1B layer of the SF3 pipe at (d) 1.5 mm, (e) 4 mm, and (f) 7 mm from the substrate. ... 104 Figure 5.38 - Boride morphologies of the SM-14Cr-1B layer of the SF3 pipe at (a)
1 mm, (b) 2 mm, (c) 4 mm, and (d) 7.5 mm. (Deep etching) ... 105 Figure 5.39 - OM images showing the prior austenite grain size of the SMSS layer
of the SF3 pipe at (a) 16 mm, (b) 22 mm, and (c) 30 mm from the substrate. ... 106 Figure 5.40 - Grain size profiles of the SF2 and SF3 pipes. ... 107 Figure 5.41 - Schematic representation of (a) the relationship between the
surface temperature profile and (b) the equilibrium temperature profile of spray-formed pipes. ... 108 Figure 5.42 - Pseudo-binary phase diagram of the SM-1B showing the equilibrium
liquid fraction at two different equilibrium temperatures... 109 Figure 5.43 - Schematic representation of the proposed model for grain size
evolution in spray forming. ... 110 Figure 5.44 - Schematic representation of (a) the surface temperature and (b)
equilibrium temperature profiles without overlapping between the SM-14Cr-1B and SMSS spray cones. ... 113 Figure 5.45 - Schematic representation of (a) the surface temperature and (b)
equilibrium temperature profiles with overlapping between the SM-14Cr-1B and SMSS spray cones. ... 114 Figure 5.46 - Schematic representation of the gradient chemical composition of
the gradient zone for an arbitrary overlapping between the SM-14Cr-1B and SMSS spray cones. ... 115 Figure 5.47 - (a) Pseudo-binary phase diagram of the 0.03C13Cr-5.8Ni-2Mo-XB
Figure 5.48 - Effect of the decreasing boron content within the gradient zone on the alloy's solidus and liquidus temperature, and the ideal equilibrium temperature profile to produce a low porosity level interface. ... 117 Figure 5.49 - Dilatometry curve of the spray-formed SMSS layer. Heating rate: 5
ºC/min. ... 120 Figure 5.50 - Amount of equilibrium phases as function of temperature showing
the presence of χ-phase in temperature range of 720-800 ºC. ... 120 Figure 5.51 - OM images showing the microstructure of the (a) as-spray formed
and (b) quenched SMSS layer. ... 121 Figure 5.52 - X-ray diffraction patterns of the as-spray formed and quenched
SMSS layer. ... 121 Figure 5.53 - Rockwell C Hardness of the as-spray formed and quenched SMSS
layer. ... 122 Figure 5.54 - Tempering curves directly from the as-spray formed conditions of
(a) SMSS, and (b) SM-14Cr-1B layer. ... 123 Figure 5.55 - OM images showing the microstructure of the SMSS layer tempered
at (a) 630 ºC for 2 h, (c) 650 ºC for 2h, and (e) 650 ºC for 6h; and SM-14Cr-1B layer tempered at (b) 630 ºC for 2 h, (d) 650 ºC for 2h, and (f) 650 ºC for 6h. ... 124 Figure 5.56 - Tensile test results of the SMSS layer (a) as-spray formed and (b)
tempered at 650 ºC for 2h. ... 126 Figure 5.57 - (a) Photography and (b) SEM image of the cup-and-cone fracture
(specimen 4) of the as-spray formed SMSS. (c) and (d) dimpled fracture surface. ... 127 Figure 5.58 - (a) Photography and (b) SEM image of the 45 degrees fracture
(specimen 1) of the as-spray formed SMSS. (c) Shrinkage solidification pore and (d) dimpled fracture surface. ... 128 Figure 5.59 - (a) Photography and (b) SEM image of the cup-and-cone fracture
Figure 5.60 - (a) Photography and (b) SEM image of the 45 degrees fracture (specimen 4) of the as-spray formed SMSS. (c) and (d) Shrinkage solidification pore. ... 130 Figure 5.61 - Charpy-V impact resistance at room temperature of the as-spray
formed and tempered SMSS layer. ... 131 Figure 5.62 - (a) Photography of the Charpy-V fracture surface and (b), (c) and
(d) mixed cleavage/dimple fracture surface of the as-spray formed SMSS. ... 131 Figure 5.63 - (a) Photography of the Charpy-V fracture surface and (b) dimpled
fracture surface of the spray-formed SMSS after tempering at 650 ºC/2h. (c) and (d) Shrinkage solidification pores. ... 132 Figure 5.64 - Tensile test results of the SM-14Cr-1B layer (a) as-spray formed
and (b) tempered at 650 ºC for 2h. ... 134 Figure 5.65 - (a) Photography and (b)-(c) SEM image of the brittle fracture
(specimen 1) of the as-spray formed SM-14Cr-1B. (d) Longitudinal view of the fracture. ... 136 Figure 5.66 - (a) Photography and (b)-(c) SEM image of the brittle fracture
(specimen 1) of the SM-14Cr-1B after tempering at 650 ºC for 2 h. (d) Longitudinal view of the fracture. ... 136 Figure 5.67 - Impact resistance of the as-spray formed and tempered SM-14Cr-1B layer. ... 137 Figure 5.68 - (a) Photography and (b) SEM image of the fracture surface of the
as-spray formed SM-14Cr-B after impact test. ... 138 Figure 5.69 - (a) Photography and (b) SEM image of the fracture surface of the
tempered SM-14Cr-1B after impact test. ... 138
1 INTRODUCTION
The extraction of oil in the pre-salt layers in the coast of Brazil brought new challenges concerning the development of materials that can withstand the severe work conditions (mainly wear and corrosion issues) found during drilling and exploitation operations. Drilling risers (steel pipes that links the well head on the sea floor to the platform) and casings (steel pipes used to protect the wall of wells during the drilling operations), for example, are frequently subjected to severe wear conditions due to the contact with rotating tool joints of the drill pipes, as represented in Figure 1.1. Moreover, these pipes work in contact with rock debris conduct by the drilling fluids (water or oil based fluids containing mainly bentonite), which increases the wear damage. Such drilling fluids are usually rich in chlorides, increasing the susceptibility of failure caused by wear and corrosion. Figure 1.2 shows examples of failed drilling risers and casings subjected to severe wear and corrosion conditions.
Figure 1.2 - (a) Wear grooves caused by the contact of rotating tool joints in the inner wall of a casing. (b) Drilling riser failed by thickness loss caused by wear and corrosion.
and M3B2 (where M = Fe, Cr, Ni, Mo) alters significantly the chemical composition
of the steel’s matrixes, resulting in changes of the stables phases of the systems. For instance, the boron-modified superduplex stainless steel presented a matrix preferentially austenitic and, high boron contents in the ferritic stainless steel resulted in a final martensitic microstructure. Secondly, the question of how the spray-formed alloys could be applied as pipe coatings was brought to light.
2 LITERATURE
2.1 Spray Forming and Co-Spray Forming Fundamentals
Spray forming is an advanced casting process in which the molten metal is directly converted to an homogeneous solid with refined structure. As shown schematically in Figure 2.1, spray forming comprises two steps: (i) atomization: the melt stream is gas atomized to produce a spray of 10 to 500 μm-diameter alloy droplets. Under action of the atomizing gas the droplets are accelerate up to 100 m.s-1 and cooled during the flight at typically 102 to 104 K.s-1; and (ii)
deposition: fully liquid, partially solid and complete solid droplets are deposited onto a substrate generating a growing spray-formed deposit. During the deposition the cooling rate of the spray-formed deposit is relatively lower, typically from 0.1 to 10 K.s-1, depending on the alloy properties and the process conditions
applied [8,9]. Billets, sheets, rings and tubes can be successfully spray formed by using appropriate substrate geometry and relative movements, see Figure 2.1 [10–21].
the equilibrium temperature were presented and only microstructures with considerable grain coarsening after solidification were shown. Recently, H. Henein [43] argued that solidification must occur in discrete regions and that there cannot be any liquid merging between adjacent droplets at the deposition zone. The argument is because the same eutectic fraction (lower than the equilibrium fraction) was observed in both the impulse atomized droplets and in the deposit formed by impulse spray of the Al-0.61%wt.Fe alloy. The lower eutectic fraction suggests that eutectic undercooling is taking place. The author suggests that when the droplets are atomized, they are covered by a nano-thick oxide coating which is not broken when the droplets impact the deposition zone, preserving what the author calls the “droplet region”. According to the author, solidification at the deposition zone of the “droplet regions” continues independently of the solute in adjacent “droplet regions”. When the deposit cools further, each “droplet region” must nucleate its own second phase, achieving the same fraction of eutectic as the atomized droplets. Nevertheless, the formation of the nano-thick oxide layer was not validated. Moreover, this model does not explain the formation of the equiaxed grains characteristic of spray-formed alloys.
a comparison between the thermal gradient of spray-formed billets in different stages of the deposition process for two different alloys: CuSn6 and 100Cr6 tool steel (AISI 52100). The numerical simulations were performed using approximately the same process conditions (see Table 2.1) and the differences in the thermal history arises mainly from the different materials properties [10,50].
Table 2.1 - Materials properties and standard boundary conditions used for the numerical simulation of spray forming of CuSn6 and 100Cr6 steel billets [10,50].
CuSn6 100Cr6 (AISI 52100)
Liquidus temperature [K] 1325 1724
Solidus temperature [K] 1189 1570
Latent heat of solidification
[kJ.kg-1] 200 287
Average thermal conductivity
[W.m-1.K-1] 153 30
Density [kg.m-3] 8484 7810
Average specific heat [J.kg
-1.K-1] (Cu, 1023 K-1301 K) 478
640 (1570 K-1800 K)
724 (1570 K-1400 K) Average liquid fraction of the
spray 0.5 0.5
Average temperature of the
impinging spray [K] 1295 1648
Convective heat transfer coefficient (billet surface )
during spray [W.m-2.K-1] 280
ℎ𝑔
= ℎ𝑚𝑎𝑥exp[−1.65 (DD max) + 0.85 (DD
max) 2
]
where D= distance to the top surface of the billet and Dmax=
reference distance = 400mm
Convective heat transfer coefficient (billet surface )
after spray [W.m-2.K-1] 10 10
Temperature of ambient air
and spray chamber [K] 523 523
Emissivity of the billet surface 0.18 0.5 Coefficient heat transfer
between billet and substrate
[W.m-2.K-1] 1000 1000
Initial temperature of the
Figure 2.2 - Numerical simulation results showing the overall temperature at different times of (a) CuSn6 and (b) 100Cr6 steel (AISI 52100) billets spray-formed using similar process conditions. Total time of the spraying process: 360 s [10,50].
for CuSn6 and 100Cr6 steel, respectively. The high thermal conductivity of the cooper alloy allows easy homogenization of the billet temperature leading to uniform solidification, that is essential for obtaining an uniform microstructure. On the other hand, after the spraying period (>360 s) the billet cools down slowly due to the high temperature gradient caused by the low thermal conductivity of 100Cr6 steel. In this case, the residual liquid is enclosed by the completely solidified material and, if shrinkage is suppressed, residual stress may rise and initiate hot cracks.
Whereas the materials properties are fixed, the process parameters can be varied in order to optimize the thermal history of the spray-formed deposits and, consequently, their final metallurgical quality. Three major parameters can be changed: (i) the specific enthalpy of the impacting spray; (ii) the heating loss during the spray process; and (iii) the heating loss after the spraying period. The specific enthalpy of the spray can be easily changed by varying the melt superheat and/or the gas to metal ratio (GMR), which results in change of the average liquid fraction of the impacting spray. Heat loss during the spray process is influenced by the temperature of the ambient gas and the initial substrate temperature. Preheating systems are often used to control the substrate temperature and influences the heat flux at the bottom of the deposit in the initial stages of the deposition process. Preheating systems are mainly important for production of spray-formed plates, strips and tubes, which are relatively thin and the region near the substrate cannot be neglected. During the cooling after the spraying period the convective heat transfer and the environment temperature can be easily controlled. For instance, by maintaining the gas flux on the deposit surface after the end of the melt flux, the convective heat transfer at the surface of the deposit can be considerably increased, and the thermal gradient and cooling rate of the deposit considered altered.
of 2.5 Hz and translates to the left at a speed of 2 mm.s-1, collecting the impinging
Figure 2.3 - Surface temperature distribution and temperature distribution through the longitudinal section of tubular preform of 100Cr6 steel (AISI 52100) under standard spray condition at spraying time of: (a) and (d) 30 s; (b) and (e) 90 s; (c) and (f) 120 s [20].
Finding the process parameters that result in the best spray-formed product quality can be challenging. Hardly the set of parameters such as, atomization pressure, GMR, atomization distance, etc. can be extrapolated from one equipment to another. This because different spray-forming equipment, with different atomizer and chamber designs, will produce different mass flux and heat flux distribution for the same set of process parameters. However, V. Uhlenwinkel and N. Ellendt [45] have shown that the surface temperature of the deposit during the deposition process is an important value to predict the porosity level of the spray-formed product. Regardless the processing parameters, the authors proposed that for a specific material, the same porosity level is achieved if the deposit surface temperature is kept constant. Figure 2.5 presents an example of porosity level as function of deposition surface temperature from collected data of several spray forming runs of Ni superalloy rings. The porosity values vary with the deposit surface temperature by a V-shaped behavior, with higher porosity values found when low deposit surface temperatures are measured. Meyer et al. [16] have shown that the same relationship between the porosity level and the deposit surface temperature can be observed in spray-formed sheets and different materials such as Cu alloys (Al-bronze and Sn-bronze) and nitriding steel. Moreover, by using the concept of dimensionless enthalpy of deposit surface (h*surf), the authors have shown that the lower porosity levels are always
Figure 2.5 - Porosity versus deposit surface temperature of spray-formed Ni superalloy rings [45].
Figure 2.6 - Effect of the dimensionless enthalpy of deposit surface on porosity of Al-bronze, Sn-bronze and nitriding steel [16].
alloys by free fall atomizer. When the substrate, tube or plate-shaped, moves through the two sprays successively, a two-layer deposit in the corresponding shape is obtained. Such multi-layer deposits are of great interest in applications where different set of properties is necessary in a single product [51]. When the two sprays are separated and parallel to each other, as represented in Figure 2.5, both spray cones do not interact to each other before they reach the substrate. In this case, the interface between the two deposited layers is relatively flat with an abrupt transition from one layer to another. Figure 2.8 presents a co-spray formed tube, which combines a Co-based alloy (Stellite 21) and a hot working steel (AISI H13) to be used in a blade for the hot forming industry. The abrupt transition between both layers can be clearly seen and a porous interface is observed by liquid penetrant inspection [53]. The major challenge in the co-spray forming process is to create an interface between the two different alloys without a porous zone.
Figure 2.8 - Cross sections of a Stellite21/AISI H13 steel co-spray formed tube processed with different parameters and the results of their liquid penetrant inspection [53].
Figure 2.9 - Schematic of co-spray forming of a gradient deposit from two different alloys using scanning gas atomizers [51].
2.2 Supermartensitic Stainless Steel
Since the early 1980s, the use of 13%Cr martensitic stainless steel grades was widely accepted in OCTG (Oil Country Tubular Goods) segment, such as pipelines, casings and risers, because of their excellent corrosion resistance, mainly in oil and gas wells containing some level of carbon dioxide [54]. However, the use of conventional type of 12-13% Cr steels, such as AISI 410 and AISI 420, has the inconvenience of their limited weldability, which requires preheating prior to welding and post weld heat treatment (PWHT) [55]. In 1990s, the duplex and superduplex stainless steel grades were extensively applied in pipelines replacing the conventional martensitic steel grades [54]. In the same period the weldable 13%Cr martensitic stainless steel grades, also called “Super 13Cr” or supermartensitic stainless steel (SMSS), were developed with improved resistance to general and localized corrosion and to sulfide stress cracking (SSC) [55,56]. Because of its enhanced properties and price, considerably cheaper than the duplex and superduplex grades, the SMSSs became an economical choice for combating CO2 corrosion and mildly sour conditions [54,57]. Between 1996
seamless pipes, 4.4% centrifugally-cast pipes and 4.3% produced by laser or other welding process [54,57].
The SMSS grades are based on the Fe-Cr-Ni-Mo system with up to 13 %wt. of Cr, 2-7 %wt. of Ni, 0.1-2.5 %wt. of Mo, low amounts of carbon, nitrogen, phosphorus and sulfur (C ≤ 0.03 %wt., N, P, S ≤ 0.03 wt.%) [55,56]. The main metallurgical concept of this steel is to increase the effective Cr content by reducing carbon to ultra-low contents (<0.03%wt.), which diminishes considerably the precipitation of M23C6 type carbides (where the main constituent
of M has been reported to be Cr). Since the low carbon content reduces considerably the hardenability of the steels, the possibility to improve the weldability by restraining the hardness of the heat-affected zone (HAZ) is also expected. However, the reduction of carbon content should be accompanied to the addition of Ni, a strong austenite-stabilizing element, in order to maintain the martensitic phase without δ-ferrite [55]. Moreover, molybdenum content up to 2.5%wt. is added aiming to improve the general, localized and sulfide stress cracking corrosion resistance.
SMSSs are normally divided into three types: (i) Lean-alloyed grade (11Cr2Ni); (ii) Medium-alloyed grade (12Cr4.5Ni1.5Mo); and (iii) High-alloyed grade (12Cr6Ni2.5Mo) [56]. Table 2.2 outlines the typical chemical composition for the three grades. Beyond the good weldability using industrial welding techniques, such steel grades have been designed to meet requirements in respect of: (i) stress corrosion cracking resistance in the presence of H2S; (ii)
diagram presented in Figure 2.11 shows that the presence of Mo narrows remarkably the region of martensitic single phase.
Table 2.2 - Typical chemical composition of SMSS grades [56]. 11Cr2Ni (lean) 12Cr4.5Ni1.5Mo (medium) 12Cr6.5Ni2.5Mo (high)
C (max %wt.) 0.015 0.015 0.015
Mn (max %wt.) 2.0 2.0 2.0
P (max %wt.) 0.030 0.030 0.030
S (max %wt.) 0.002 0.002 0.002
Si (max %wt.) 0.4 0.4 0.4
Cu (max %wt.) 0.2-0.6 0.2-0.6 0.2-0.6
Ni (%wt.) 1.5-2.5 4.0-5.0 6.0-7.0
Cr (%wt.) 10.5-11.5 11.0-13.0 11.0-13.0
Mo (%wt.) 0.1 1.0-2.0 2.0-3.0
N (max %wt.) 0.012 0.012 0.012
Table 2.3 - Alloy design of the SMSS grades to meet target corrosion resistance [57]. Environmental parameter 11Cr2Ni (lean) 12Cr4.5Ni1.5Mo (medium) 12Cr6.5Ni2.5Mo (high)
Temperature 20-100 ºC 20-100 ºC 20-100 ºC
P (CO2) 10 bar 20 bar 20 bar
P(H2S) - 0.005 bar 0.050 bar
pH 3.5-4.5 3.5-4.5 3.5-4.5
Figure 2.10 - (a) Effect of molybdenum on CO2 corrosion resistance at 180 ºC for
two different nickel concentration. (b) Effect of molybdenum on the resistance to sulfide stress-corrosion cracking (applied stress: 100% yield strength). Base composition 0.025C-13Cr-0.45Mn (wt%.) [58].
Figure 2.11 - Experimental diagram showing the boundaries of the austenite, ferrite and martensite phases as function of Cr, Ni and Mo concentration for 0.01 wt.% C after austenitization at 1050 ºC and air cooling [55].
the other hand, the presence of retained austenite is reported to be beneficial to the impact toughness by means of a localized transformation-induced plasticity (TRIP) effect [60–62]. However, it has been shown that the presence of high fraction of retained austenite increases the hydrogen solubility of the alloy, which increases the susceptibility of hydrogen-induced failure [63,64]. In sub sea oil and gas pipelines, the main hydrogen sources are cathodic protection and hydrogen in weld metal. Figure 2.12 shows the impact resistance at low temperatures (chary-V tests) of rolled plates of the three typical SMSS grades after quenching and tempering heat treatments. One can see that when the microstructure is optimized trough heat treatment and no δ-ferrite is present, the three grades of SMSS present remarkable impact resistances at sub-zero temperatures. Table 2.4 shows the typical tensile properties of the three SMSS grade after such optimized heat treatment [56]. Some authors reported that the mechanical properties of the SMSS grades can be increased by addition of low amounts of Ti and Nb (about 0.1 %wt.) due to the precipitation of fine titanium/niobium carbonitrite particles in the martensitic matrix [65,66]. Moreover, adding Ti/Nb decreases the amount of Cr-rich precipitates, as both combine preferentially with residual nitrogen and carbon to form nano-scale precipitates during tempering, which is also beneficial to the corrosion resistance.
Table 2.4 - Mechanical properties of the three SMSS grades. Hot rolled plates after optimized heat treatments with tempering temperatures close to AC1 [56].
11Cr2Ni (lean) 12Cr4.5Ni1.5Mo (medium) 12Cr6.5Ni2.5Mo (high) Yield Stress
Rp0.2 (MPa) 641 738 705
Tensile Stress
(MPa) 826 889 878
5d Elongation
Figure 2.12 - Impact resistance (Charpy-V) at sub-zero temperatures of the three typical SMSS grades. Optimized microstructure free of δ-ferrite. The hot rolled plates were heated over 800 ºC and water quenched. The optimum balance between strength and toughness is achieved after tempering in a narrow range of temperatures close to AC1. Typical grain size is ASTM: 7-10.
2.3 Boron modified stainless steels
In spite of good mechanical and corrosion properties, stainless steel grades are usually known for presenting low wear resistance. Several applications such as tubing, pumps and valves found in oil exploitation and production require materials that combine high mechanical and corrosion resistance to high wear resistance. Typical examples are the drilling casings and risers which must present: i) high mechanical properties to withstand the high stresses; ii) high corrosion resistance due to the marine environment and the often presence of CO2 and H2S contents; and iii) high wear resistance due to the contact of the
rotating drilling pipe to its inner wall. However, it is a big challenge to find a material presenting this set of properties. Based on this, many developments of alloys to be applied as wear and corrosion resistant coatings to conventional steels have been performed.
the research group of the Materials Engineering Department at the Federal University of São Carlos has reported several developments of boron modified stainless steels processed by spray forming [1,4,5,7]. The addition of boron to the chemical composition of stainless steel grades leads to formation of hard borides, mainly M2B-type borides (where M is composed of the transition metal
usually present in stainless steels such as Fe, Cr, Ni, Mo), while maintaining the features of the stainless steel matrix. In 2011, Beraldo, L. [1] reported the first developments on boron modified superduplex stainless steel (SDSS). Figure 2.13 shows the microstructure of the SDSS modified with 3.7 %wt. of boron processed by spray forming. The addition of high boron content led to formation of high fraction of primary M2B-type borides embedded in the duplex
ferrite/austenite matrix with some fraction of Mo-rich eutectic borides (M3B2-type).
However, changing in the chemical composition by boron addition resulted in a much higher volume fraction of austenite in the matrix, as high as 90%, far from the usual 50% of conventional SDSS grades. Although such results have shown that controlling the microstructure of boron-modified SDSS may be difficult, the presence of the hard M2B-type borides in the stainless steel microstructure yields
an abrasion wear resistance, measured by the rubber wheel against dry sand method (ASTM G65), as good as Stellite 1016 coatings (commercial wear resistant Co-based alloy). Such promising result led to a series of further developments on different boron-modified stainless steel grades.
Nascimento, A.R.C [4] has reported the effect of different boron contents (1.0, 2.0 and 4.0 %wt.) on the microstructure and wear resistance of spray-formed ferritic AISI 430 stainless steel. The author has shown that changing the boron content in the alloy can considerably alter their solidification path resulting in completely different microstructures. When 1.0 and 2.0 %wt. of boron are added, the alloys solidification comprises the formation of primary δ-ferrite, followed by the peritectic reaction δ + L →M2B. During the peritectic reaction, Cr diffuses from
the edges of the primary dendrites toward the borides yielding a final microstructure composed of borides embedded in ferritic matrix with a Cr-poor zone around the borides (see Figure 2.14 (a) and (b)). On the other hand, the addition of 4.0 %wt. of boron leads to the formation of primary M2B borides, which
reduces the Cr content of the remaining liquid. When boron is completely consumed by the boride formation, the final remaining liquid has a composition within the γ-austenite field where the solidification is completed. Due to the high levels of alloying elements and the cooling rate of the small spray-formed deposit, the final alloy microstructure comprises large primary borides embedded in a martensitic matrix. The results presented by Nascimento, A.R.C. [4] showed that addition of high boron contents makes more difficult the maintenance of the stainless steel matrix characteristic of the base alloy, which may reduce the corrosion properties as, for instance, in the case of the alloys presenting Cr-poor regions. However, despite the differences between the final microstructures, the wear resistance (evaluated by a plate-on-cylinder wear test) of the boron modified AISI 430 stainless steels are considerably higher than the low alloy high strength steel API X80 (often used in risers and casings manufacture) [4].
Aiming at the achievement of a boron-modified stainless steel grade with high wear resistance and the maintenance of the features and corrosion properties of the base alloy, the development of the boron-modified supermartensitic stainless steel with lower boron contents (up to 1.0 %wt.) was proposed [5]. The first results showed that addition of small boron contents in the SMSS medium alloy grade processed by spray forming results in a microstructure composed of equiaxed martensitic grains with eutectic M2B-type borides at the grain boundaries (see
Figure 2.15). Moreover, when the boron content is increased two effects can be observed: (i) increase in the borides fraction; and (ii) reduction of the equiaxed grain size. Both effects have shown to be beneficial to the alloys' abrasive wear resistance [5,7] . In this doctoral thesis, the complete designing of the spray-formed boron-modified SMSS is present. The microstructural evolution, the relationship between the boride fraction and the wear resistance in different wear mechanisms, and the effect of boron addition on the corrosion properties will be addressed.
3 OBJECTIVES
4 MATERIALS AND METHODS
In order to achieve the stated objective, this work was divided into three parts with well defined goals. Firstly, the comprehensive development of the spray-formed boron-modified SMSS grade will be presented aiming: (i) to understand the relationship between the solidification features of the spray forming process and the final microstructure of the boron-modified SMSS; (ii) to address the relationship between the microstructure of spray-formed boron-modified SMSS grades and their wear resistance; and (iii) to address the effect of the boron addition on the corrosion resistance of the spray-formed boron-modified SMSS with different chemical compositions. Secondly, bimetallic pipes of boron-modified SMSS (inner layer) and conventional medium-alloyed SMSS (outer layer) were produced in a semi-industrial spray-forming plant aiming at: (i) verifying the effects of the larger scale process on the final microstructure of the boron-modified SMSS and the conventional SMSS; and (ii) evaluating the possibility to achieve a high quality product in terms of porosity and microstructural homogeneity. Finally, the mechanical properties of both layers of a bimetallic pipe obtained were evaluated in the as-spray-formed condition, and after heat treatments, aiming to evaluate which level of mechanical properties is possible to be achieved directly from the spray-forming process, which means without further thermomechanical treatments that is beyond the scope of this thesis. The description of the materials and experimental methods applied to each part is described in the next sections.
4.1 Design of Spray-formed boron-modified SMSS 4.1.1 Laboratory Scale Spray Forming
iron-boron alloy with 16%wt. of iron-boron, iron-molybdenum alloy with 62 %wt. Mo, commercial pure chromium and nickel were used as raw materials. Table 4.1 shows the aimed chemical composition of the four alloys. In the 0.3B, SM-0.7B and SM-1B, the selected composition aimed at adding boron while maintaining the same Cr, Ni and Mo content of the commercial medium-alloyed grade of the SMSS used as base alloy. On the other hand, in the SM-14Cr-1B composition, the addition of 1 %wt. of boron was accompanied by an increase of the Cr content up to 14 %wt. while keeping the same levels of Ni and Mo.
In each spray forming run, approximately 4 kg of raw materials were melted in an induction furnace and spray formed onto a rotating carbon steel disc substrate using N2 as atomization gas. The pouring temperature of all materials
was 1650 ºC, the spray distance 460 mm and the substrate speed 45 rpm. The average melt flow and gas flow of all spray forming runs were approximately 0.133 kg.s-1 and 0.170 kg.s-1, respectively, resulting in a gas-to-metal ratio (GMR)
of approximately 1.2.
Table 4.1 - Aimed chemical composition of the four spray-formed boron-modified SMSS.
SM-0.3B
%C %Cr %Ni %Mo %B %Si %Mn %S
0.01 11.90 5.80 2.00 0.3 0.25 0.45 <0.01
%P %W %Co %Cu %V %Nb Ti %Fe
0.02 0.01 0.03 0.06 0.03 0.01 0.13 Bal.
SM-0.7B
%C %Cr %Ni %Mo %B %Si %Mn %S
0.02 11.90 5.80 2.00 0.7 0.25 0.45 <0.01
%P %W %Co %Cu %V %Nb Ti %Fe
0.02 0.01 0.03 0.06 0.03 0.01 0.13 Bal.
SM-1B
%C %Cr %Ni %Mo %B %Si %Mn %S
0.02 11.90 5.80 2.00 1.0 0.25 0.45 <0.01
%P %W %Co %Cu %V %Nb Ti %Fe
0.02 0.01 0.03 0.06 0.03 0.01 0.13 Bal.
SM-14Cr-1B
%C %Cr %Ni %Mo %B %Si %Mn %S
0.03 14.00 5.80 2.00 1.00 0.25 0.45 <0.01
%P %W %Co %Cu %V %Nb Ti %Fe
0.02 0.01 0.03 0.06 0.03 0.01 0.12 Bal.
4.1.2 Microstructure Characterization
The chemical compositions of the final alloys were determined by inductively coupled plasma atomic emission spectrometry-ICP-AES, excepting C and S that were analyzed by direct combustion. Phase’s identification was performed by XRD analysis using a Rigaku Geigerflex ME210GF2 model diffractometer with Cu-Kα radiation. The microstructures were characterized by scanning electron microscopy (SEM) using a FEI Inspect S50 Scanning Electron Microscope. In order to reveal the microstructures the polished samples were etched with a 3HCl:1HNO3 solution. Deep etching with 10 mL HCl, 3 ml HNO3, 5
mL FeCl3 and 82 mL ethyl alcohol solution were performed in order to reveal the
4.1.3 Equilibrium Solidification path determination
Thermodynamic simulations by using the Thermo-Calc software were performed to determine the equilibrium solidification path of the boron-modified SMSS grades. The database used was the TCFE7. The pseudo-binary phase diagram of the composition Fe-12%Cr-5%Ni-2%Mo-X%B (%wt.) was calculated using the equilibrium calculator function. The same function was used to calculate the equilibrium phase fractions at different temperatures, ranging from 1000 ºC to 1600 ºC, of the boron-modified SMSSs with the exact chemical composition of the spray-formed deposits produced.
It is worth stressing that recent works of our research group (not published yet) have shown that thermodynamic simulations using Thermo-Calc is an useful tool to determine the equilibrium solidification path of boron-modified stainless steels. However, it is important to point out that such works have shown that since the thermodynamic data available is usually obtained for steels containing low boron contents (< 0.1 %wt. of boron), the solidus and liquidus temperatures of iron-based alloys with higher boron contents (> 1%wt.) differ considerably from the experimental values.
4.1.4 Wear tests
Wear test specimens were machined from equivalent radial positions of the spray-formed deposits. Two different wear testing were performed in order to evaluate the wear resistance of the spray-formed boron-modified SMSSs. Firstly, the dry sand against rubber wheel abrasive wear test was performed in accordance with the procedure A of ASTM G65-04 standard.
during the wear tests this initial normal force is slightly increased due to changing in the relative position of the mass center of the arms as consequence of the thickness loss of the sample. This increase of the normal force can be accurately described as a function of the loss thickness (h) by the following polynomial equation: FN = 539.34 + 7.1267h + 0.0865h2 + 0.0004h3, where the units of FN
and h are N and mm, respectively; and these increments were considered to perform the calculus of contact pressure. In all tests, the rotation speed used was 252 rpm. In order to simulate the wear conditions found in risers and casing, the chambers were filled with 6 liters of drilling fluid donated by System Mud ltda. with the composition shown in Table 4.2. Three samples of each material were tested and the values presented here are the mean value and standard deviation of these samples.
Figure 4.2 - Schematic representation of the plate-on-cylinder wear test.
Table 4.2 - Composition of the drilling MUD used in the plate-on-cylinder wear tests.
Content Bentonite
KCl
Viscosifier polymer Sand (AFS 50/70)
When the cylinder slides in contact with the sample, a crescent worn groove is formed as shown in Figure 4.2. The groove volume, or the accumulated worn volume, was measured at each 30 min along 10 hours and plotted as function of the sliding distance. The obtained curves were fitted using equation 4.1, which is similar to the empirical model presented by Hall and Malloy [69] to describe the wear behavior of casings and risers in real scale wear tests.
𝑉 = 𝐴{1 − 𝑒𝑥𝑝[−𝐵(𝑠
𝑐)]}(4.1)
Where V is accumulated worn volume in m³, s is the sliding distance in m and A, B and C are constants parameters that describe the evolution of wear with the sliding distance. As reported in [69], this function in equation 4.1 can represent a remarkable variety of shapes, all of which have one characteristic in common: As the sliding distance increases, the accumulated worn volume approaches to the limiting value A. In other words, this function shows that the wear rate decreases with the increasing of the sliding distance. Physically, reaching the limit A value does not mean that the wear completely stops, but that the wear rate reached very low values, which can be neglected. After fitting the experimental data with equation 4.1, the wear rate (dV/ds) can be calculated by using the equation 4.2.
𝑑𝑉
𝑑𝑠
= 𝐴. 𝐵. 𝐶. exp[−𝐵(𝑠
𝑐)] . 𝑠
(𝑐−1)(4.2)
According to the authors [69], the decrease of the wear rate is associated with the reduction of the contact pressure caused by the increase of the contact area when the groove grows. When the limiting A value is reached the contact pressure also reach a constant value, so called threshold contact pressure (TCP). The TCP is now an important wear property, and the lower the TCP values obtained and the faster they are reached in a system, the better is the materials wear resistance.
Figure 4.3 - Microstructure of the commercial SMSS bar used as reference alloy for the wear tests. Etching: Vilela.
4.1.5 Electrochemical corrosion tests
Electrochemical impedance spectroscopy (EIS) and potentiodynamic polarization analyses were performed using a conventional three electrodes set up. As counter-electrode (CE) and reference electrode (RE) a platinum sheet and a saturated calomel electrode (SCE) were used, respectively. A schematic representation and a photography of the electrochemical cell used are presented in Figure 4.4. The analyses were carried out using a Gamry potentiostat. A solution of 35 g/L of NaCl in deionized water and pH = 4.0 (controlled by addition of H2SO4) was used as electrolyte. Three test protocols were applied:
(I) The samples were immersed 24 h in deionized water for a pre-passivation treatment. After pre-passivation, it was carried out 1 hour of open circuit potential (OCP) measurements in the electrolyte. Subsequently, EIS analysis, with potential amplitude of 10 mV around the OCP value and frequencies varying from 105 Hz to 10-2 Hz, was performed. Since EIS is a non-destructive technique, after
EIS analysis, the sample was kept 10 minutes longer in OCP, and potentiodynamic polarization was carried out in sequence. The potentiodynamic polarization curves were obtained by sweeping the potential from 50 mV below the corrosion potential to a maximum potential (named critical potential - Ecrit)
corresponding to a current of 0.1 mA.cm-2.
(II) The samples were directly subjected to 1 hour of OCP measurements, without pre-passivation treatment. Subsequently, the EIS and polarization tests were carried out following the same protocol described above.
(III) The samples were maintained 12 hours in solution while measuring the open circuit potential (OCP). Subsequently, the EIS and polarization tests were carried out following the same protocol described in protocol (I). In this case, after EIS measurements the samples were left 1 hour in OCP before the polarization test.
To observe the corroded surfaces of the spray-formed boron-modified SMSSs, the samples with polished surfaces (using alumina suspension with particle size up to 1 µm) were polarized anodically until the current density has reached 5 mA.cm-2. The corroded surfaces were observed by optical microscopy
Figure 4.4 - (a) Schematic representation and (b) photography of the electrochemical cell used for potentiodynamic and EIS analyses.
Electrochemical noise (EN) analyses were carried out in order to evaluate the repassivation kinetics of the spray-formed boron-modified SMSS after a simulated wear event, in the same line of the scratch tests that have been used since several years [70]. Two identical samples of the same alloy were connected as CE and WE at the electrochemical cell in a zero resistance ammeter, ZRA, configuration. After assembling the electrochemical cell, both samples were maintained for 24 hours immersed in the solution. The coupling current between both samples was measured during the last hour of immersion before being interrupted. The surface of the sample acting as CE was then abraded through a #240 grit sandpaper for 10 seconds (using a hand drill with constant rotation), simulating an abrasive or wear event upon the metallic surface. Immediately after finishing the abrasion, the coupling current record was resumed to the survey of the repassivation tendency.
Table 4.3 - Chemical composition of the commercial supermartensitic stainless steel (%wt.).
%C %Cr %Ni %Mo %Si %S %P %Mn
0.007 11.86 5.87 2.00 0.25 0.001 0.016 0.45
%Nb %W %Co %N %Cu %V Ti %Fe
0.01 0.01 0.03 0.0095 0.06 0.03 0.135 Bal.
4.2 Spray forming of semi-industrial scale bimetallic pipes 4.2.1 Spray Forming runs
Three spray forming runs hereinafter named SF1, SF2, and SF3 were carried out at the SK1+ spray forming plant in the IWT-University of Bremen. The Sk1+ spray forming plant has two sets of induction heating furnaces (with capacities of 5 L and 12 L) and pouring systems, which are used to delivery and atomize simultaneously different alloys by free fall atomizer. Figure 4.5 shows the schematic representation of the SK1+ spray forming set up. A rotating cylindrical preform was used as substrate. The boron-modified SMSS and the medium-alloyed SMSS are atomized simultaneously while the substrate translates in such a way that the boron-modified is firstly deposited. The SMSS is then deposited onto the previously deposited boron-modified layer forming the bimetallic pipe.
In the SF1 run, AISI 316L stainless steel was used as feedstock material. On the other hand, commercial medium-alloyed SMSS was used as base material in the SF2 and SF3 runs. Fe-Mo alloy, Fe-B alloy, Fe, Ni and Cr (commercially pure grades) were added in order to obtain the aimed chemical composition. Table 4.4 shows the chemical composition of the feedstock materials. The amount of material used in each run and the target chemical composition of each spray-formed alloy are shown in Table 4.5 and 4.6, respectively.
Table 4.4 - Chemical composition (%wt.) of the feedstock materials.
AISI 316 L
%C %Si %Mn %Cr %Ni %S %P %Mo
0.03 0.402 1.78 16.00 10.00 0.031 0.039 2.00
%Nb %W %Co %Cu %V %B Ti %Fe
0.004 0.048 0.163 0.336 0.056 0 0.011 69.11
SMSS
%C %Cr %Ni %Mo %Si %S %P %Mn
0.007 11.86 5.87 2.00 0.25 0.001 0.016 0.45
%Nb %W %Co %N %Cu %V Ti %Fe
0.01 0.01 0.03 0.0095 0.06 0.03 0.135 Bal.
Fe*
%C %Si %Mn %Cr %Ni %S %P %Mo
0.005 0.05 0.2 0.2 0.3 0.025 0.025 - %Cu %Fe
0.2 98.995
Fe-Mo %C %Si %S %P %Mo %Cu %Fe
0.028 2.23 0.0288 0.1 62.68 0.34 34.59 Fe-B %C 0.326 0.57 %Si %S 0.002 %P 0.034 %B 16.54 82.528 %Fe
Cr** %Cr 100
** Commercially pure iron from Höganäs (Compacted powder)
Table 4.5 - Amounts of feedstock material used in each spray-forming run.
SF1 SF2 SF3
SM13Cr1B SMSS SM14Cr1B SMSS SM14Cr1B
AISI 316L (kg) 12.945 29.956 - - -
SMSS (kg) - - 20.414 64.100 20.414
Fe (kg) 10.812 28.671 - - -
Cr (kg) 1.4289 3.274 0.930 - 0.930
Ni (kg) - - 0.150 - 0.150
Fe-Mo (kg) 0.178 0.389 0.066 - 0.066
Fe-B (kg) 1.634 - 1.439 - 1.439
Total weight (kg) 27.00 62.90 23.00 64.10 23.00
Table 4.6 - Target chemical composition of the spray-formed alloys.
SF1 SF2 SF3
%wt. SM13Cr1B SMSS SM14Cr1B SM14Cr1B SMSS
C 0.034 0.017 0.027 0.027 0.01
Cr 13.00 13.00 14.5 14.5 11.86
Ni 4.90 4.95 5.87 5.87 5.84
Mo 1.37 1.35 19.95 19.95 2.00
B 1.00 - 1.00 1.00 -
Fe/minor
elements Bal. Bal. Bal. Bal. Bal.
tilted 12º in in order to reduce the distance between both spray cones (as shown in Figure 4.5). In addition, the atomizer of the SMSS was scanned ± 6º relatively to the axial axis of the cone increasing the range of the second spray cone. The substrate was not preheated and its initial temperature was the room temperature. In this run, it was not possible to measure the surface temperature of the deposit during the process.
Based on the results obtained from the SF1, the SF2 run was set up aiming to improve the quality of boron-modified SMSS layer (as it will be discussed further in the results). In the SF2 run, only the first layer of boron-modified SMSS was spray-formed. The main difference from the SF1 to the SF2 run is the utilization of a preheat system as depicted in Figure 4.5. In the SF2 run, the substrate was preheated at 950 ºC before starting the process. In order to counterbalance the hotter process conditions imposed by the substrate preheating, the metl flow rate was decreased by reducing the nozzle diameter from 5.0 mm to 4.0 mm, which resulted in higher GMR (about 3.2).
Finally, in the SF3 run, both layers were spray-formed again. The substrate was preheated at 950 ºC and the process parameters applied to the atomization of the boron-modified SMSS was similar to those applied in SF2. To counterbalance the hotter process conditions imposed by the substrate preheating, the parameters applied to the SMSS layer were also changed. Both the atomizer gas pressure and the nozzle diameter were reduced from 0.6 MPa and 6.5 mm to 0.5 MPa and 6.25 mm, respectively. In addition, in this experiment, the atomizer was not scanned. Such set of parameters resulted in higher GMR (0.79) when compared to the SF1 run (0.58).
In the SF2 and SF3 runs the surface temperature of the deposit was measured during the process by an optical pyrometer positioned in the region of the SM-14Cr-1B cone as shown in Figure 4.6
